Full-Scale Seismic Testing of Piles in Improved and Unimproved Soft ClayFebruary, 2016
Full-Scale Seismic Testing of Piles in Improved and Unimproved Soft Clay Bradley Fleming, Iowa State University Sri Sritharan, Iowa State University Gerald A. Miller, University of Oklahoma Kanthasamy K. Muraleetharan, University of Oklahoma
Available at: https://works.bepress.com/sri_sritharan/34/
Bradley J. Fleming,a) M.EERI, Sri Sritharan,b) M.EERI, Gerald A. Miller,c)
and Kanthasamy K. Muraleetharanc)
A full-scale field investigation was performed to determine the effects of soil improvement on the seismic resistance of piles in soft clay. A soil improvement method, called cement deep soil mixing (CDSM), was used to improve soil supporting a standard 324 mm diameter steel pipe pile subjected to simulated earthquake lateral loads. An identical pile in unimproved clay was also tested to determine the effects of the soil improvement. Compared to the unimproved pile, the CDSM technique showed a 42% increase in pile lateral strength, a 600% increase in effective elastic stiffness, and a 650% increase in average equivalent damping ratio. The pile in improved soil reached its lateral capacity at a head displacement of 0.1 m, at which point the critical region at the base of the pile above the improved ground experienced buckling and subsequent fracture due to low cycle fatigue. [DOI: 10.1193/012714EQS018M]
INTRODUCTION
Structures and their foundations are subjected to forces created by earthquakes, wind, waves, water current, vessel impact, ice, and gravity. All loads applied to the superstructure must be transmitted to the foundation, which are then transferred to the surrounding soil. With significant lateral loads, such as those created by earthquakes, use of pile foundations is one of a few options to transmit large structural loads to competent soils. Piles supported by competent soils are relatively easy to design cost-effectively. However, thick layers of weak soils such as soft clays are widespread in high seismic areas (e.g., San Francisco, southern Nevada, Washington, Eastern Missouri, and Arkansas), exacerbating design challenges. Driven piles are the preferred supports for many structures found on saturated soft soils due to ease of installation and their effectiveness at penetrating deep layers of competent soils as needed. However, soft clays reduce the lateral resistance of the pile-soil system, making the pile foundation less cost-effective. In this case, the current design practice is to use an increased number of larger diameter piles (see SDC, Caltrans 2010), which may not be feasible in deep deposits of soft clay. An innovative and more cost-efficient solution to this problem is to improve the soil surrounding the pile in the immediate vicinity
a) Department of Civil, Construction, & Environmental Engineering, Iowa State University, 394 Town Engineering, Ames, IA 50011
b) Department of Civil, Construction, & Environmental Engineering, Iowa State University, 351 Town Engineering, Ames, IA 50011
c) School of Civil Engineering and Environmental Science, University of Oklahoma, 202 West Boyd Street, Norman, OK 73019
Earthquake Spectra, Volume 32, No. 1, pages 239–265, February 2016; © 2016, Earthquake Engineering Research Institute 239
over a short depth, thereby increasing its lateral stiffness and allowing the pile lateral strength to be fully developed.
Some well-known methods for improving the soil conditions in the field include deep soil mixing, jet grouting, stone columns, and simple soft soil replacement. However, soil improvement techniques are not often used in design practices due to limited understanding of the behavior of improved soil and interactions between the pile, improved soil, and unim- proved soil. In addition, this limited understanding often results in overly conservative soil improvement designs that include large improvement volumes and higher cost as compared to an efficient design. Most experimental and analytical studies focused on utilizing soil improvement to mitigate liquefaction of loosely deposited sand but without the presence of piles (Mitchell et al. 1998, Martin et al. 2001, Hatanaka et al. 1987, Adalier 1996, Adalier et al. 1998, Iai et al. 1988, Akiyoshi et al. 1993, Liu and Dobry 1997, Kawakami 1996). Seismic behavior of piles in liquefiable sands has also been extensively studied (e.g., Ashford et al. 2000a, Ashford et al. 2000b, Weaver et al. 2005, Boulanger and Tokimatsu 2006, Ohtomo 1996), while a majority of studies for piles in soft clay have been investigated by the offshore community (Vucetic and Dobry 1988, Basack and Purkayastha 2007). Only a few studies have addressed the seismic behavior of pile foundations constructed on soft clays (Brown et al. 2001, Wilson 1998, Boulanger et al. 1999, Meymand 1998, Lok 1999, Mayoral et al. 2005), despite the widespread presence of the soft clay soil in high seismic regions and the frequent need to locate bridges and buildings in this soil type. Only recently, a few investigations have been carried out to determine the effectiveness of ground improvement on increasing the lateral resistance of pile foundations embedded in soft clay.
Rollins and Brown (2011) improved the quasi-static and dynamic behavior of full-scale pile groups in the field by treating soft clay in the vicinity of the pile and/or the pile cap using several different soil improvement techniques including jet grouting, soil mixing, flowable fill, soil replacement, and rammed aggregate piers. Generally, soil improvement methods using cement as the stabilizing agent (e.g., jet grouting, soil mixing, and flowable fill) pro- duced the largest increases in lateral resistance compared with other soil treatment methods. The jet grouting technique was applied to the soil region surrounding the piles below the pile cap in a block type configuration. Neglecting the passive resistance of the pile cap against the soil, the lateral resistance of the system increased by 220% compared to an unimproved pile group. In addition, the initial stiffness of the system increased by a factor of about 14 based on the results from small displacement tests (<0.03m).
These experimental studies show that improvement methods using cement to treat the soil are viable techniques for significantly increasing the resistance of piles in soft soil. However, limited experimental research has been performed to isolate the effects of soil improvement surrounding single piles and to demonstrate the impact that soil improvement has on enhan- cing the seismic behavior of piles in soft clay. To address this issue, a full-scale experiment using dynamic and quasi-static lateral loads applied at the pile head was conducted on two piles having identical pile properties and site conditions but one was embedded in cement deep soil mixing (CDSM) improved soil and the other was not. Responses of the improved pile are computed in terms of percent increases in key response characteristics as compared to the unimproved pile to demonstrate the improved seismic behavior of CDSM treated piles in
240 FLEMING ET AL.
soft clay. To assist in estimating the response of CDSM improved piles, recommendations are given to estimate the static responses of the improved soil based on back-calculation of experimental data. The method is verified using a computer program called LPILE (Ensoft 2010) to compare computed soil responses to the observed responses. Finally, the impact of soil improvement is demonstrated using LPILE to determine the wall thickness and diameter of an unimproved pile to achieve the same response of a standard pile in improved soil that is taken from the experimental results.
TEST SITE
For the field experiment, a soft soil site in Miami, Oklahoma, was chosen. The test site consisted of a 4.4 m layer of soft clay overlying a 2.0 m layer of sandy gravel and limestone bedrock. The top 1.0 m consists of lean clay with gravel and occasional construction debris. The soft clay at the site was classified as lean clay (CL) according to the Unified Soil Clas- sification system (ASTM D2487). Soil characterization at the site included 11 piezocone soundings (CPTu), two soil borings, Shelby tube sampling, and a stand pipe piezometer to measure the depth of the water table (Taghavi et al. 2010).
Figure 1 shows the geotechnical characteristics of the soil profile at the test site. Undrained shear strength (su) and effective friction angle ( 0) of the soil were calculated from CPTu tip resistance (qt) and pore-water-pressure (U) using Equations 1 and 2, respec- tively (NCHRP Synthesis 368):
0 50 100 150 200
S u (CPTu)
S u (Triaxial)
S u (kPa)
q t
q t (MPa)
Lean clay with gravel and occasional debris
Medium stiff to very soft silty clay (30 LL, 12 PI)
Sandy gravel
Ground surface
Figure 1. Average geotechnical characteristics of the Miami, Oklahoma, test site.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 241
EQ-TARGET;temp:intralink-;e1;41;640su ¼ qt σvo Nkt
(1)
qtσatmffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi σvo
0σatm p
(2)
where σvo is equal to the total overburden pressure, Nkt is equal to the dimensionless cone factor, and σatm is equal to the atmospheric pressure. An average unit weight (γ) of 19 kNm3
for the clay and 20 kNm3 for the sand was used to calculate the overburden pressure. The water table, measured by a stand pipe installed at the site, was 2.62 m below the ground surface at the time of pile testing. Piezocone soundings were performed when the water table was about 0.80 m below the ground surface, as shown in Figure 1. This depth was used to determine the effective overburden pressure (σvo
0). Nkt was determined to be about 17 based on the average undrained shear strength of soft clay samples taken from the field and tested in the lab. Based on the CPTu measurements, the average undrained shear strength of the soft clay was generally in the range of 30 to 95 kPa between the depths of 1.1 m and 4.4 m below the ground surface. Several Shelby tube samples were tested in the lab for consolidated undrained shear strength of the soft clay soil that are shown in Figure 1 and are detailed in the site investigation report (Taghavi et al. 2010). Generally, the shear strengths produced in these tests were between 35 kPa and 60 kPa for depths between 1.5 m and 2.6 m. Also determined in the lab were the overconsolidation ratio (OCR), the liquid limit (LL), and the plastic index (PI) for the clay, which are also shown in Figure 1. At depths greater than 4.4 m, the average friction angle of granular soil was between 40 degrees and 45 degrees based on CPTu soundings.
TEST UNIT DETAILS AND CONSTRUCTION
A three-dimensional (3-D) rendering of the full-scale test configuration is shown in Figure 2. Two identical steel pipe piles having AISC size HSS12.75 0.375 and satisfying ASTM A106B specifications were installed in improved and unimproved soft clay. The
Improved Soil
Reaction Frame
Soft Clay
(a) (b) Reaction Frame
soil not shown
Figure 2. (a) Soil profile and test configuration and (b) picture of test site during dynamic testing of pile in improved soil.
242 FLEMING ET AL.
average Young’s modulus and yield strength for the steel in each test pile, determined from four tensile tests performed on samples cut from the sidewall of each pile, were 213 GPa and 372 MPa, respectively. A yield plateau in the stress-strain curve initiated at an average strain of 0.17% and terminated at an average strain of 1.1%, at which point the steel exhibited strain hardening. The average ultimate strength of the steel was 588 MPa, which occurred at an average strain of 15%. The total length of each pile was 7.6 m. Approximately 5.2 m of each pile was embedded below the ground surface and into the underlying sandy gravel layer to help support the pile and reduce the amount of full length rotation under lateral loading. A hole was dug around each test pile, near the ground surface, having approximate dimen- sions of 1.2 m in diameter and 1.0 m in depth, to remove the effects of the clay with gravel layer and to assist in improving the uniformity of the clay in contact with the pile.
A pile cap, consisting of two halves of a concrete block and acting as a seismic mass needed for dynamic testing, was clamped to the head of the pile to be tested. A set of threaded rods, sent through the pile cap, avoiding the pile in the center, were used to clamp the pile cap to the pile and to support a stiffened C15 40 channel used for connecting the quasi-static and dynamic actuators. The mass was kept constant for all tests to isolate the effect of varying natural frequency of the system. Although observing changes in the natural frequency of the system would benefit determining the effects of superstructure mass, the volume of concrete necessary to achieve the effective mass of a bridge deck or building would be unfeasibly large to include in the experiment. Therefore, a smaller mass was included to achieve a target natural frequency within the testing range so that reasonable estimates of system damping, which are best found at the natural frequency of the system, could be determined.
Construction of the improved soil volumes occurred before the installation of piles. The improved soil was constructed using CDSM. In this process, an ABI Mobilram machine with an augur drive attachment was used to revolve a mixing tool into the soft clay. It consisted of a hollow shaft and mixing paddles, as shown in Figure 3a. Cement grout was pumped through the hollow shaft and ejected laterally behind the lower mixing paddle, where it was mixed with the native soil. While still revolving, the mixing tool was advanced to a depth of 4.0 m and retracted. This process was repeated once more to form a well-mixed column of soil and cement. The top of the improved soil column after mixing is pictured
(a) (b)
Figure 3. Pictures of (a) Mobilram and auger with mixing paddles and (b) top of soil improve- ment column.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 243
in Figure 3b. Concrete mixer trucks transported the grout from the batch plant to the site and dumped the grout into a hopper where the grout was continuously mixed and pumped through the supply line for the mixer. Water from a nearby river was also pumped into the hopper periodically to maintain cement and water concentration close to the target value.
The improved volume was conservatively designed to test the full strength of an im- proved pile relative to an unimproved pile. The objective was to improve the soil such that the test pile would yield and form a plastic hinge as opposed to rotating under the lateral load. The nominal dimensions of the improved volume were 13D 13D 13D, where D is equal to the outer diameter of the test pile. Because of the presence of the hole around the base of the pile, a distance of approximately 9D along the pile length denotes the contact length between the pile and improved soil.
To construct the large improvement block, a single column was constructed, after which the mixing tool was offset slightly to achieve an overlap of approximately 0.3 m between columns. The process was repeated to form a block of improved soil with the CDSM column arrangement shown in Figure 4. A total of 16 columns with diameters of 1.2 m were con- structed, resulting in a plan dimension of 4.0 4.0m. A test pile was then installed in the center of the improved volume while the improved soil was still wet using the same ABI Mobilram machine but with a vibrating hammer attachment. This process allowed the pile to be installed with ease.
In the case of retrofitting an existing foundation with CDSM, although not as effective as the block type configuration chosen for the experiment as demonstrated by Rollins and Brown (2010), the improved soil could be constructed as a wall in close proximity to the foundation to increase soil resistance to lateral loading of the foundation elements includ- ing piles and pile cap. If possible, jet grouting, which also uses cement grout to improve soil resistance, can be applied to the soil underneath the pile cap and adjacent to piles via holes drilled through the cap to allow access to this region. These retrofit configurations were
3962 mm
3962 mm
Test Pile
Figure 4. Mobilram (left) and CDSM column arrangement and test pile placement (right).
244 FLEMING ET AL.
studied by Rollins and Brown (2010) for pile groups and resulted in significant increases in system stiffness compared to an unimproved pile group. As a basis for future development of CDSM improved piles, the purpose of the current research is to isolate the effects of soil improvement surrounding a single pile in absence of an underground pile cap.
The grout used to construct the CDSM columns consisted of a water-to-Portland-cement (Type I) ratio of 11 by weight. A total of 1.893m3 of cement grout was added to each column, resulting in an approximate concentration of 20% cement by dry weight of soil in the CDSM column. The average undrained shear strength of the improved soil, determined in the lab using samples collected from the field, was 1517 kPa. This was computed using a compressive-strength-to-shear-strength conversion factor of 3.2 as opposed to the conven- tional value of 2, typically used for clay. The revised value developed from testing of clay soil with high cement concentrations in direct shear (Porbaha et al. 2000). These authors estab- lished the correlation shown in Equation 3 as a result of their research, which demonstrated a quadratic relationship between the shear strength (τf 0) and compressive strength (qu) in units of kgfcm2 for the improved soil. However, it is important to note that loading conditions for simple shear are difficult to achieve in the laboratory and that results from low normal stress conditions used in the tests are more appropriate for shallow soil-cement:
EQ-TARGET;temp:intralink-;e3;62;425τf 0 ¼ 0.53þ 0.37qu 0.0014q2u for qu < 60 kgfcm2 (3)
Figure 5 shows the plan and profile views of the test specimens and reaction frame. The reaction piles, having AISC size HP10 42 and conforming to ASTM structural grade A572-50, were installed outside the zone of influence of the test pile in unimproved soil (TPU) and the test pile in improved soil (TPI) that, according to the Canadian Geotechnical Society (1978), was taken as a radial distance of 10D away from the center of the test pile. The two reaction piles closest to TPI were driven into two separate col- umns of improved soil to increase the strength of the reaction frame. Approximately 6.2 m of each pile was embedded below the ground surface and into the underlying sandy gravel layer.
A steel frame, consisting of HP10 42 steel members and mounted to the four reaction piles as shown in Figure 5, was designed to support the dynamic actuator while maintaining an appropriate distance between the actuator and the pile cap. A removable cantilever seg- ment, which provided a platform for mounting the dynamic actuator to the frame, was designed to support the dynamic actuator on both sides of the main reaction frame for testing of both piles. After removing the cantilever segment, the actuator for quasi-static testing was mounted to a cross beam supported by the reaction piles of the frame (shown on the right side of Figure 2a). All members and connections within the reaction frame were designed to support a maximum 888 kN lateral force, after considering a safety factor of 2, which was applied at the center of the clevis for the dynamic actuator just above the can- tilever and at the center of the quasi-static reaction beam, separately. Frame size and con- figuration was based on soil properties determined from multiple CPTu soundings at the site. The shear strength of the soil was reduced by 50%, which conservatively accounts for the variability of soil strength in the field and resulted in a conservative reaction frame design.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 245
INSTRUMENTATION
Figure 5 shows the locations of instrumentation installed at the site. Each test pile was instrumented with over 32 strain gages, four displacement transducers, three tilt meters, and three accelerometers. Steel angles, having a size L2 2 18, were welded to the test piles to safeguard strain gages and cables during driving. Generally, tack welds were placed every 305 mm along the length of the test pile, in between the strain gage locations to reduce the risk of damaging them. Displacements of each test pile were measured at three different elevations along the freestanding portion of the pile. A fourth displacement transducer was placed at the same elevation as the top transducer but with a horizontal separation of 300 mm to measure possible twisting of the pile cap during testing. All displacement transducers were mounted to a reference beam placed adjacent to the test pile and supported
5.26 m
Accelerometers
1.83 m
(b) Profile View
Figure 5. Plan and profile views of full-scale test configuration.
246 FLEMING ET AL.
by hollow structural members, which were driven outside the zone of influence, again, taken as a distance of 10D away from the pile (Canadian Geotechnical Society 1978). Forces applied to the pile cap were measured using the load cells integrated into the actuators. Three accelerometers located on the ground surface, which were placed adjacent to the test piles, were used to measure the vibration transmitted from the pile to the soil. Subsurface accelerometers were also installed but only near the unimproved pile. Pore-water-pressure transducers were also installed below ground near the unimproved pile.
LOADING PROTOCOL
Dynamic and quasi-static testing protocols were selected for each pile to evaluate the cyclic behavior of the test piles at varying levels of loading. Generally, test piles were sub- jected to at least three cycles of the same amplitude and frequency, after which the amplitude was increased and the next load pattern was applied. Figure 6 shows the loading protocol used for both test piles. TPU was first subjected to dynamic loading with force control enabled. Displacement magnitudes were between 0.2mm and 115mm. Initially, fre- quency of the excitation was varied between 0.25 Hz and 8 Hz for displacement magnitudes of less than 2mm. However, due to the velocity limitations of the actuator, excitation frequencies were reduced with increasing displacement magnitudes, as noted in Figure 6. For instance, displacement magnitudes of 50mm were achieved with 2 Hz excitation fre- quency while the displacement cycles at 115mm were applied at a frequency of 0.4 Hz. Undergoing this process also helped with understanding the pile response as a function of frequency.
Quasi-static loading was performed to test piles at larger lateral displacements. Displace- ment magnitudes less than the maximum displacement reached during dynamic testing were applied to TPU in the early stages of quasi-static testing. Only one cycle of the load pattern was applied for each interval of increasing displacement up to the maximum displacement of 115mm, after which three cycles of symmetric loading were applied at each interval as the displacement magnitude increased to a maximum of 406mm. This was the last symmetric load pattern, which was dictated by the stroke limitations of the actuator. The remaining cycles displaced the pile 406 mm in the push direction and 566 mm in the pull direction from center.
Initially, TPI was subjected to small vibrations with force control enabled. Force control was limited to magnitudes smaller than 2.5 kN. Subsequently, displacement control was enabled for displacement magnitudes between1mm and100mm. Frequency of the exci- tation was varied between 0.25 Hz and 8 Hz for displacement magnitudes of 1mm. For displacement magnitudes larger than 1mm, a load pattern sequence was applied to TPI with similar increasing displacement magnitude increments and excitation frequencies as TPU. The target quasi-static load pattern for TPI was also similar to the load pattern chosen for TPU. However, the pile failed before completing the entire load pattern sequence.
TEST OBSERVATIONS
Figure 7 shows some of the key test observations. Under cyclic lateral loading, soft clay surrounding TPU experienced significant plastic deformations, evident by soil losing contact with the pile (i.e., gapping) after the pile head was brought back to zero displacement.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 247
100 10
1 0.1
0.01 0.1
1 10
Cumulative Full-Cycle Sine Waves
0 2 4 6 8 10 12 14 16 21 23 26 28 41 44 47 50 53 74
D is
p la
ce m
en t
(m m
10 F
re q
u en
H z)
Additional cycles of same magnitude and frequency not shown for clarity
Force Control
Displacement Control
Cumulative Full-Cycle Sine Waves
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20
D is
p la
ce m
en t
(m m
10
Cumulative Full-Cycle Sine Waves
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19
D is
p la
ce m
en t
(m m
*
(d)
Figure 6. Test sequences used in following load protocol: (a) TPU during dynamic testing, (b) TPU during quasi-static testing, (c) TPI during dynamic testing, and (d) TPI during quasi-static testing.
Conversely, the improved soil surrounding TPI remained relatively intact during all tests. Less than 3 mm total measured gap width occurring at the ground surface and short hairline cracking, which ran parallel to the direction of loading, was observed in the improved soil directly in front and behind the pile. TPI reached its lateral capacity at a displacement of about 10 cm, at which point the critical region of the pile about 75 mm above the surface of the improved ground experienced buckling and fracture due to low cycle fatigue.
GAP OPENING
The soil surrounding TPU had undergone significant plastic deformations, causing gap- ping between the pile and the soil during unloading. Following each critical load cycle, the width of gap was measured manually using a scale at the ground surface after the pile was brought back to zero displacement. Figure 8 shows average gap width versus average pile head displacement for multiple cycles of varying head displacements of TPU, suggesting a linear trend. This was also observed in Suleiman et al. (2006) during lateral load tests of a drilled shaft in frozen and another drilled shaft in unfrozen soil. A linear regression through
(a) (b) (c)Hairline Crack
Figure 7. Test observations including (a) gapping of soil adjacent to TPU, (b) hairline cracks in improved soil adjacent to TPI, and (c) fracture of TPI after 10 cm displacement.
Pile Head Displacement (mm)
G ap
W id
th (m
y = 0.54x - 11
Figure 8. Measured gap width at the ground surface for TPU.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 249
the data shows a y-intercept of −11mm, which suggests the soil adjacent to the sidewall of the pile at the ground surface generally displaced 11 mm in the same direction as the pile immediately after reversing the load. The rebounding effect of the soil occurs regardless of displacement magnitude. This was verified by comparing the displacement of the pile near the ground surface and the measured gap width, which, on average, was a difference of about 11 mm. This evidence is also present in the back-calculated responses of soil displacement. From these measurements, the soil rebound could be observed in higher resolution. It was found that soil rebound increased with increasing displacement of the pile up to a pile dis- placement magnitude of about 37 mm, after which the rebound remained relatively constant at 11 mm. Gapping for TPI was negligibly less than 3 mm for all quasi-static tests.
FORCE-DISPLACEMENT RESPONSES
The graphs in Figures 9 show the dynamic and quasi-static force-displacement responses obtained at the pile head for TPU and TPI. These responses show distinctively different hysteresis behaviors, which are largely caused by the introduction of improved soil in TPI. In addition, varying excitation frequency had negligible effect on the global force- displacement response of the system, as can be seen in the response envelope, as no abrupt change to the stiffness or force resistance is seen under increasing excitation frequency. Equivalent viscous damping ratios (ξeq) were calculated using Equation 4 (Chopra 1995) for each cycle at varying pile head displacements to quantify the inelastic energy dissipa- tion (i.e., system damping with the linearly elastic damping component removed) in both systems:
EQ-TARGET;temp:intralink-;e4;41;367ξeq ¼ 1
-200
-150
-100
-50
0
50
100
150
200
250
-250
-200
-150
-100
-50
0
50
100
150
200
250
Py - = -152 kN(a) (b)
Figure 9. Global force-displacement response of (a) TPU and (b) TPI.
250 FLEMING ET AL.
where the inelastic energy dissipation is equal to the area within the force-displacement hys- teresis loop (ED) of the quasi-static data and ES is equal to the maximum strain energy in the system or ksecu2m2 with a system stiffness ksec and maximum displacement um. Figure 10 shows ξeq for quasi-static tests in TPI and TPU except for tests associated with the þ400mm 475mm cycle of TPU due to the asymmetric loading. Global responses for each system are detailed in the paragraphs below.
UNIMPROVED PILE (TPU)
The strength and stiffness of TPU, determined from the quasi-static response of the sys- tem, were estimated by obtaining lateral applied force and head displacement at the first occurrence of yielding in the pile, which corresponds to a positive strain of 0.17% developing along the outer surface of the pile. The pile first yielded 1.32 m below the ground surface at an applied lateral head force of 129 kN, termed herein as system strength (Py), and a lateral head displacement of 17 cm. Therefore, the secant stiffness of the system was 759 kNm. At the beginning of the experiment, TPU was loaded with equal displacements in the push and pull directions of loading. However, the midpoint of the full extension and retraction range of the actuator was about 9 cm closer to the reaction frame compared to the centerline of the pile, which caused the force-displacement plot of the last three cycles to be unsymmetrical in full range of the actuator displacement.
Generally, ξeq increased with increasing head displacement except for lateral head dis- placements less than 11 cm, where previous dynamic testing had disturbed the surrounding soil. In this region, ξeq decreased with increasing head displacement because ksec of the sys- tem increased as lateral head displacement approached the previous maximum head
Figure 10. Equivalent viscous damping ratio for independent hysteresis loops during quasi-static testing.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 251
displacement of 11 cm, which subsequently increased the elastic strain energy stored in the system. In addition, ξeq decreased with increasing number of cycles for any given head dis- placement. This is attributed to an increase in gap width between the pile and soil with increasing number of cycles and hence less soil is engaged during repeated cycles.
IMPROVED PILE (TPI)
The first occurrence of yielding in TPI was at a negative head displacement (pull) of 2.6 cm, which is an 85% decrease in yield displacement compared to TPU. Correspondingly, Py was 152 kN in the pull direction, which is 18% higher than Py for TPU, which is a result of the plastic hinge located 1.0 m closer to the ground surface. The stiffness of TPI, estimated using the same method to find the stiffness of TPU, was 5;846 kNm, which is over seven times greater than the effective stiffness of TPU. The pile also yielded in the push direction, which was at a lateral head displacement of about 3.6 cm. The force applied at the top of the pile was 191 kN, resulting in an effective stiffness of 5;306 kNm, taken from the origin of the force-displacement curve to the yield point.
Compared to TPU, ξeq is larger for TPI and is more consistent with an increasing number of cycles, which suggests that almost all energy dissipation was through the plastic action of the pile. For head displacements less than 5 cm, ξeq values were less than 8%. On occasion, ED computed from testing TPU was larger compared to ED values of TPI, which was observed in the areas within hysteresis loops, especially at larger displacements. Although energy dissipation was larger in the system for TPU, so was ES due to large lateral displace- ments of the pile head, resulting in smaller ξeq values by comparison.
RESPONSE PROFILES
Figure 11 shows the strain profiles for selected head displacements in the push direction for TPU and TPI. Plastic zones developed in TPU and TPI at the locations corresponding to strains larger than the yield strain (εy), which is equal to 0.17%. For simplicity, only the strain gage measurements on the extreme tension side of the pile when the pile is pushed are dis- cussed in this section. Generally, the strain profiles on the extreme compression side in the push direction mirror the measured strains shown in Figure 11 when the pile remains elastic. However, strains larger than εy are influenced by the loading history and the cyclic moment- curvature response of the pile section. This is the reason for the observed increases and decreases in strain measurements on the extreme tension side of TPU and TPI for increasing head displacement beyond the limit required to yield the pile. However, the strain measure- ments on the extreme compression side (not shown) change appropriately to follow the moment-curvature behavior of the section. Important parameters including plastic hinge loca- tion and length can be observed in the data provided and are noted for each system in the following paragraphs.
Maximum strain in TPU occurred 1.95 m below the ground surface and the plastic zone near this location was approximately 2.5 m long when a lateral head displacement of 400 mm was applied in the push direction. Strains were also sufficiently large to cause a plastic zone to occur along a short segment near the ground surface. The steel angles protecting the strain gages were terminated at this location, which caused a change in the flexural rigidity (EI) of the pile with respect to distance along the pile length. The EI of the pile was not only affected
252 FLEMING ET AL.
by the presence of the angles but also by the location and distance between welds connecting the angles to the pile. Based on the back-calculation of EI from strain gage measurements and applied lateral loads, the elastic EI was found to increase from 25;400 kN-m2 (EI for a pile cross section without angles) at the termination of the angles to a depth of about 2.2 m below the ground surface, at which point the maximum EI of 33;400 kN-m2 for a cross section including angles was achieved. This increase in EI also increased the moment required to yield the pile based on theoretical moment-curvature analyses of two cross sections. One pipe section without angles and another with angles were analyzed for flexural strength based on properties of the steel determined in the lab. Bending forces sufficient to cause a pipe section to yield without and with angles were 258 kN-m and 285 kN-m, respectively. The moment in TPU, at 10 cm below ground, was within this range based on strain gage measurements observed in Figure 11a. Therefore, the pile underwent yielding at this location. However, the rate of increase in EI with respect to depth was greater than the increase in the ratio of yield moment over the curvature in the pile. Therefore, the pile did not yield at loca- tions where the product of curvature and EI were less than the yield moment. Once EI reached its maximum value, yielding occurred again at the maximum moment location and the plastic zone increased in length from this location with increasing load.
The maximum strain in TPI was located just below the ground surface. The sidewall of TPI experienced inelastic strains at head displacements larger than 36 mm in the push direc- tion, even in phase 1 testing, which occurred during the 50mm displacement cycle, as can
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Figure 11. Strain profile of (a) TPU and (b) TPI.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 253
be seen in Figure 11b. The plastic hinge that developed in TPI was approximately 90 cm long after the 200mm displacement cycle and extended 61 cm below the ground surface. Also note that strain measurements were a constant zero below a depth of approximately 1.5 m, which is within the improvement zone. This observation suggests that the depth of the improved soil could have been reduced by about 50% along the region where the pile is in contact with the improved soil.
BACK-CALCULATING SOIL RESPONSE
Soil response can be back-calculated analytically from experimental data assuming that soil displacement (y) and soil resistance (p) at discrete layers along the length of the pile are equal to pile displacement and the load distribution along the length of the pile, respectively. Theoretically, based on Euler–Bernoulli beam theory, y and p are calculated by double inte- gration of the pile’s curvature and double differentiation of the pile’s moment, respectively, with respect to the distance along the length of the pile. If the pile behaves elastically, the moment profile can be calculated in terms of cross-sectional behavior of the pile as the pro- duct of flexural rigidity and curvature of the pile. In this study, the theoretical flexural rigidity of the pile was calculated and applied to discrete values of curvature obtained from strain gage measurements. Brandenberg et al. (2010) have shown that the accuracy of p is greatly affected by the method used to differentiate the data. These authors further found that a dif- ferentiating technique of minimizing weighted residuals gives comparable performance to the method of fitting cubic splines (Segalman et al. 1979) to the data and outperforms the curve fitting of the higher-order polynomials method (de Sousa Coutinho 2006). The weighted- residual method performed slightly better than the cubic spline method when noisy data are sampled using small intervals along the pile length. The higher-order polynomials method produced accurate derivatives for uniform soil profiles but was not suitable for layered soil. Alternatively, a generalized constraint function can be used for soil response and optimized based on pile response and cross-sectional behavior as demonstrated by Tehrani et al. (2014). However, this method requires a constraint function for soil behavior, which has not been established for improved soil. For these reasons, the weighted-residual method was used to back-calculate soils response in the current research.
WEIGHTED-RESIDUALS METHOD
Suppose there is a desired function gðzÞ equal to the derivative or integral of a function f ðzÞ with respect to the distance z along the pile length. As mentioned previously, only dis- crete values f i are known at locations of zi. Correspondingly, gi are discrete values obtained from the weighted-residuals method. Complete details of the differentiation method can be found in Brandenberg et al. (2010). The integration method can be performed in a similar manner, which has not been described in previous literature. Therefore, both methods are presented below in a compact form as
EQ-TARGET;temp:intralink-;e5;41;154g ¼ df dz
≈ ½G ¼ 3½Z1½Δ½F; (5)
EQ-TARGET;temp:intralink-;e6;41;111g ¼ ð L
3 ½Z½Δ1½F; (6)
254 FLEMING ET AL.
f n1
2 66666664
. .. . ..
. . . .
2 66666664
2ðz2 z1Þ ðz2 z1Þ 0 0 · · · ðz2 z1Þ 2ðz3 z1Þ ðz3 z2Þ 0 · · ·
0 ðz3 z2Þ 2ðz4 z2Þ ðz4 z3Þ · · ·
. . . ..
. .. . ..
. . . .
· · · 0 ðzn1 zn2Þ 2ðzn zn2Þ ðzn zn1Þ · · · 0 0 ðzn zn1Þ 2ðzn zn1Þ
3 77777775 :
Equations 5 and 6 were applied twice to moment and curvature data to calculate p and y, respectively. Appropriate boundary conditions were required to determine both rotation and displacement of the pile. For small pile head displacements, as observed during dynamic testing, the displacement and rotation were zero at the bottom of the pile based on strain profile measurements. However, for head displacement magnitudes of 100 mm and larger, the strain profiles penetrated almost to the bottom of the pile, indicating potential for full pile rotation. In this case, boundary conditions were obtained using tilt sensors and displacement transducers located near the pile head for quasi-static test results. The boundary conditions for shear in the pile and soil reactions were known at the ground surface and provided an additional discrete measurement in the system of equations. The shear at the ground surface was taken from load cell measurements in the actuator and the soil reaction along the unsup- ported length of the pile was taken as zero.
FIELD P-Y CURVES
Figure 12 shows the average normalized soil resistance (ppu) for the CDSM improved soil and soft clay soil, which were calculated from strain gage measurements taken in the field, where pu is the ultimate resistance of the soil and is dependent on the effective unit weight of the soil (γ 0s), undrained shear strength of the soil (su), depth of the soil layer (z), and diameter of the pile (d). Normalizing the soil resistance by computing ppu removes the effects of the soil properties and overburden pressure such that selected curves can be plotted together and verified with models found in the literature. An appropriate theoretical pu profile was used for each pile. Using limit equilibrium methods, Matlock (1970) proposed the pu profile for clay to be computed as
EQ-TARGET;temp:intralink-;e7;62;136pu ¼ min
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 255
The model chosen to represent the unimproved soil was derived by Matlock (1970), which is a model for soft clay. To the authors’ knowledge, the derivation of a p-y model for improved soil has not been attempted by other researchers. However, the response of CDSM soil may be similar to that of stiff clay (without free water) or weak rock, which have predicted p-y curves based on the work performed by Reese and Welch (1975) and Reese (1997), respectively. Guo (2013) developed a general-use model for sand and clay, which represents soil behavior with a linear elastic, perfectly plastic p-y curve. All four of the aforementioned models were compared to the back-calculated curves shown in Figure 12.
The back-calculated resistance of the soft soil is largely nonlinear, as shown in Figure 12a. The back-calculated curve fits Matlock’s curve well for soil displacements less than 20mm, especially in the positive displacement region of the curve. According to the figure, at maximum soil displacements (jyj > 50mm), the normalized resistance of the back- calculated curve reaches an apparent maximum value less than the maximum soil resistance of the model, which is likely due to a reduction in effective overburden stress caused by the excavation of soil around the test pile. Both the model and the back-calculated curves were normalized using the same theoretical value for pu, which was computed using Equation 7, taking z as the depth from the top of the excavation. In reality, z is approximately equal to zero at the bottom of the excavation and equal to the theoretical value at a depth where the exca- vated soil has minimal effect on the overburden pressure. This effect is difficult to determine for practical purposes and thus was not accounted for when constructing Figure 12a.
Figure 12b shows the p-y behavior of the improved soil. It is seen that the resistance of the improved soil is relatively linear for small lateral displacements (jyj < 1.5mm). An attempt
Figure 12 Normalized p-y responses for (a) unimproved soil and (b) improved soil.
256 FLEMING ET AL.
was made to match the improved soil with a model developed for stiff clay (Reese and Welch 1975), as was done by Huang (2011). However, the stiffness of the model was significantly less than the stiffness back-calculated from field experiments for soil displacements greater than 1.5mm. The model developed for weak rock (Reese 1997) has an initial linear p-y stiffness (K) dependent on the Young’s modulus of the soil (Es) and a dimensionless constant kir that was derived from experimental data. However, recommended kir values for weak rock are significantly higher than the back-calculated values found in the experiment, as can be seen in Figure 12b. Despite this difference, both the model and the back-calculated curve reach pu at about the same displacement of about þ1.5mm. Both the stiff clay and rock models show large nonlinear behavior, which was not observed in the back-calculation of p-y for improved soil. The model described by Guo (2013) is linear and fits the back- calculated curve very well. According to the author, K for this model can be estimated as
EQ-TARGET;temp:intralink-;e8;62;493K ¼
0.087þeð11.49þ50eÞ (8)
where e is the ratio of the unsupported length of the pile over the length of the pile below the ground surface, Ep ¼ EIpðπd464Þ is the Young’s modulus of an equivalent solid pile, EIp is the flexural rigidity of the pile, Es is the Young’s modulus of the soil, and ν is the Poisson’s ratio of the improved soil. Es was taken as 500 MPa, which was the average modulus of stress-strain curves produced by unconfined compression tests of the improved soil from the field. This resulted in a K equal to 984MNmm assuming a Poisson’s ratio of 0.25. The pu for both the Guo (2013) model and back-calculated curve were computed using the theoretical equation shown in Equation 7, where z was taken as the depth from the bottom of the excavation.
ANALYSIS
Figures 13 and 14 show the results of a model compared to the responses of the piles observed in the field. In this analysis method, the pile was represented as a series of nonlinear beam-column elements. These elements were connected to a bed of nonlinear springs repre- senting soil response. The springs’ resistance, p, and deflection, y, were considered to vary continuously with depth. Empirical models developed by Matlock (1970) and Reese et al. (1974) were utilized for the response of unimproved clay and sand, respectively. A theore- tical model developed by Guo (2013) was used for the improved soil. The responses of each of these models can be found above and the properties of the system can be found in the details of the experiment. Recommendations for p-y curves were based on homogeneous soil layers with constant properties horizontally to infinite distances. This was not satisfied at depths where improved soil was present. However, using the response of TPI, it is shown for soil improvement of sufficient width that the model deployed in this paper was satisfactory. A computer program called LPILE, developed by Ensoft, was used to ana- lyze the lateral load behavior of a pile using the Winkler spring bed model
Generally, there was a good comparison between the field and analysis for both systems. The global responses for TPI and TPU in the field were not symmetric in terms of the force magnitudes in the push and pull directions (i.e., the magnitude of the force in the pull direc- tion is larger than the force in the push direction), whereas the analysis is symmetric. This was
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 257
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Figure 13. Comparing lateral pile head responses in the field (solid) and LPILE (dashed) for (a) TPU and (b) TPI.
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Figure 14. Comparing response profiles for (a) TPU and (b) TPI in the field and LPILE.
258 FLEMING ET AL.
likely the result of the frame and reaction piles interacting with the test pile. Since the analysis does not account for the influence of the frame, the response of the analysis does not match the response observed in the field in the pull direction. The moment profiles of the analysis also show good agreement with the field results.
EVALUATING EQUIVALENT RESPONSES
As previously described, soil improvement significantly increases the lateral stiffness and strength of a pile in soft clay and may be a cost-effective solution as opposed to increasing pile diameter and number of piles in the foundation. The cross-sectional area of the pile (Ap) is directly related to the amount of steel in the pile and approximately correlated to the cost of the pile. Increasing the diameter of the pile also increases the cost of manufacturing and installing the pile. However, the costs of materials and construction are controlled by fluc- tuating market prices across different areas of the world, which makes a detailed cost analysis difficult and outside the scope of this paper. Therefore, the motive for this section is to describe the effectiveness of CDSM soil improvement in terms of an equivalent pile in unim- proved soil but with enhanced section properties that provide the same equivalent system strength and stiffness as a pile in CDSM improved soil but with a smaller cross section. The responses of TPI were used to compare with the responses of a theoretical pile developed in LPILE (Ensoft 2010) with the same soft soil conditions observed in TPU. Relative increases in the diameter and wall thickness of the unimproved pile were provided as a result of this comparison.
The strength of the system was taken as the force applied at the pile head required to yield the steel along the outer radius of the pile. This allowed the full strength of the pile to develop while the surrounding soil provided the resistance required to resist the applied load. System stiffness was also determined at the first occurrence of yielding in the pile and was calculated by dividing the system strength by the lateral displacement at the pile head. To find the equivalent strength and stiffness of the theoretical pile in unimproved soil equal to that of TPI, the pile diameter and wall thickness were varied in LPILE until the response of the system met the following conditions at the pile head simultaneously: (1) a maximum strain in the pile equal to 0.17%, which was εy, found in testing of the steel in the lab, (2) a lateral force applied at the pile head equal to 152 kN, which was Py observed in testing of TPI, and (3) a lateral displacement of the pile head equal to 2.6 cm, which was the pile head displacement that caused yielding in TPI.
Figure 15 shows the effect of system strength and stiffness in terms of Ap for indepen- dently varying diameter and wall thickness of the unimproved pile in LPILE. As the diameter of the unimproved pile was increased and the wall thickness was held constant, the resistance of the soil and pile increased, which in turn increased the strength of the system. Increasing the wall thickness of the pile and keeping the diameter constant also increased system strength due to the additional resistance provided by the pile at yielding, but the resistance of the soil remained constant. Therefore, increasing the diameter of the pile for a given Ap had a greater effect on system strength than increasing wall thickness due to increasing resistance provided by the soil.
System stiffness was observed to have an optimal section size denoted by the peak in the arched curves shown in Figure 15b. It was observed that increasing Ap increased the applied
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 259
force required to yield the pile. Therefore, the stiffness of the soil must decrease with increas- ing Ap due to the nonlinearity of its behavior for a given pile diameter. It was also shown that pile stiffness increased with increasing Ap. The peak in the curve of constant diameter denotes the condition where system stiffness transitions from pile stiffness dominated behavior to soil stiffness dominated behavior. Therefore, the stiffness of the theoretical unimproved pile in LPILE could not achieve the equivalent stiffness of TPI for a diameter that was less than
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Figure 16. Force-displacement behavior of TPI and LPILE analysis of a pile with a 254 cm outside diameter and 0.3 mm wall thickness.
260 FLEMING ET AL.
sufficiently wide. A diameter of at least 120 cm resulted in soil resistances large enough to achieve the equivalent stiffness provided that the wall thickness was greater than 5 mm. However, the system strength corresponding to this section size was too large compared to the strength of TPI. Therefore, the wall thickness of the pile was reduced to decrease the strength of the system. However, this also resulted in the reduction of system stiffness, which had to be compensated by increasing the diameter of the pile. Through this optimiza- tion procedure, the pile diameter and wall thickness that produced an equivalent system response to that of TPI were determined to be 254 cm and 0.3 mm, respectively.
Figure 16 shows the force-displacement response of the equivalent pile in LPILE com- pared to the measured response of TPI in the field. The stiffness and strength of both systems were comparable, as expected. However, the capacity of the equivalent pile was generally 36% higher than the measured capacity observed for TPI in the field. This was attributed to a longer plastic hinge length and larger pile diameter in the equivalent pile compared to TPI.
CONCLUSIONS
With the objective of understanding and improving the seismic behavior of piles sup- porting foundations in soft soil, a full-scale field investigation was conducted by applying dynamic and quasi-static loads to piles in improved and unimproved soft clay. Both test units had identical pile dimensions and site profiles. One of the piles was embedded in a volume of improved soil (TPI), applied using the CDSM technique, and the other was left unimproved (TPU). From the experimental results presented in this paper, the following conclusions have been drawn:
1. The secant stiffness of TPU was 759 kNm at the first occurrence of yielding in the pile, which occurred at a lateral head displacement of 17 cm. The pile had undergone significant lateral displacement before the capacity of the pile was achieved. This was due to limited resistance of soft clay and as a result reduces the effectiveness of the pile at lateral displacements less than 5 cm, which is typically taken as the limit in seismic foundation design. In addition, a gap had formed between the pile and soft clay, which caused the system to become more flexible with increasing number of cycles and magnitude.
2. The impact of soft clay was minimized with the application of CDSM and increased shear and moment in the pile of TPI, allowing the pile to achieve its capacity at more reasonable lateral displacements (i.e., <5 cm). The stiffness and strength of the sys- tem were 5;846 kNm and 152 kN, respectively, at the first occurrence of yielding in the pile, after which the pile had buckled and fractured just above the ground surface due to low cycle fatigue. In addition, negligible gapping and cracking to the CDSM improved soil was observed, even after 200mm of displacement was applied to the pile head. For displacements beyond yielding condition, the energy dissipation of TPI was mostly contained within the pile.
3. With respect to the response of TPU, the improved soil introduced the following changes to the lateral load responses of TPI:
• Increased the system strength by 42%. • Increased secant stiffness at yielding condition of the pile by a factor of
approximately 7.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 261
• Increased the equivalent damping ratio by a factor of approximately 7.5. • Shifted the locationof themaximummomentupwardsbyapproximately130cm.
4. LPILE, a computer program that utilizes the Winkler spring bed model for the non- linear response of the soil in terms of soil resistance, p, and displacement, y, per- forms satisfactorily in predicting the monotonic response of a pile in unimproved soil and improved soil of sufficient width and depth. The p-y curves utilized in LPILE for unimproved soil also compare to the p-y relationship back-calculated from pile strain measurements in the field. The p-y response of the improved soil may be estimated using a linear elastic, perfectly plastic model with a stiffness (K) of 984MNmm.
5. Two criteria were met for achieving the same strength and stiffness of TPI using larger section sizes for a pile in unimproved soil. The results showed that a pile section having an outside diameter of 254 cm and a wall thickness of 0.3 mm can achieve a strength and stiffness similar to that of an improved pile. Installing a pile with this section size is not possible with the construction equipment avail- able for pile driving. However, these results do show the diameter of the pile required to engage enough soft clay to resist the same loads applied to TPI. Therefore, applying CDSM is an effective means to achieve 100% of the strength of the pile.
ACKNOWLEDGMENTS
This material presented in this paper is based upon work supported by the National Science Foundation under NEESR-SG Grant No. CMMI-0830328. Dr. Richard J. Fragaszy serves as the program manager for this grant. Any opinions, findings, and conclusions or recommendations expressed in this paper are those of the authors and do not necessarily reflect the views of the National Science Foundation.
The Oklahoma Department of Transportation, the City of Miami, and the Grand River Dam Authority provided support in locating and evaluating the soil at the test site. Juan Baez, president of Advanced Geosolutions Inc., and Arul Arulmoli, principal of Earth Mechanics Inc., provided valuable expertise and resources for the construction of the improved soil. Charbel Khoury from the University of Oklahoma and Owen Steffens and Douglas Wood from Iowa State University provided valuable assistance in constructing the test setup. Robert Nigbor, Steve Keowen, and Alberto Salamanca of nees@UCLA provided valu- able expertise and resources instrumenting and testing the piles in the field.
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(Received 27 January 2014; accepted 24 November 2014)
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT CLAY 265
February, 2016
Full-Scale Seismic Testing of Piles in Improved and Unimproved Soft Clay
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